Measurement of toughness in the heat affected zone of welded structural steels

8/19/2019 Measurement of toughness in the heat affected zone of welded structural steels

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Commission of the European Communit ies

t ech n ica l s tee l research

Proper t ies and serv ice per fo rmance

MEASUREMENT OF TOUGHNESS IN THE

HEAT-AFFECTED ZONE OF WELDED

STRUCTURAL STEELS

R e p o r t

EUR 9297 EN

Blow-up from microf iche original


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Commission of the European Communit ies

t e c hn i c a l s te e l re s e a rc h

Proper t ies and serv ice per fo rmance



STRUCTURAL STEELS

M.J. GEORGE

BRITISH STEEL CORPORATION

9 , A l b e r t E m b a n k m e n t

G B - L O N D O N S E 1 7 S N

C o n t r a c t N o 7 2 1 0 - K A / 8 0 4

1 . 7 . 1 9 7 8 - 3 1 . 1 2 . 1 9 8 2 )

FINAL REPORT

Di rec to ra te -Genera l

Sc ience , Research and Deve lopment

1985 EUR 9 2 9 7 EN


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P u b l i s h e d b y t h e

C O M M I S S I O N O F T H E E U R O PE A N C O M M U N I T I E S

D i r e c t o r a t e - G e n e r a l

I n f o r m a t i o n M a r k e t a n d I n n o v a t i o n

L - 2 9 2 0 L U X E M B O U R G

L E GAL NOT ICE

Neither the Commiss ion o f the European Communi t ies nor any person act ing

on behal f o f the Commiss ion is respons ib le for the use which might be made of

the fo l low ing in fo rm a t ion

>ECSC-EEC-Euratom, Brus se ls · L uxem bourg


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STRUCTURAL STEELS

F I N A L R E P O R T

Agreement No. 7210.KA/804

M.J. George

British Steel Corporation

Teesside Laboratories

EUR 9297 EN


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FR 62-10 822 7210.KA/804


MEASUREM ENT OF TOUGHNESS IN THE HEAT AFFECTED ZONE

OF WELDED STRUCT URAL STEELS

ECSC Agreement No. 7210.KA/804

SUMMARY

In view of the increasing importanc e of heat affected zone (HAZ) pro pert ies in

the specification and in-service behaviour of steels for demanding structural

appl icat ions , a study has been carried out, with the aim of:-

(a) Examin ing critica lly the metho ds current ly availab le

for investigating the fracture toughness of

HAZ's,

with particular reference to the assessment of fitness

for purpose and the ability to predict the potential

service performance of HAZ's from the results of small

scale te sts, by comparison with wide plate tests

containing HAZ notches.

(b) Asses sing the prop erti es attai nable in the HAZ' s of some

structural steels currently used in applications where

high HAZ property levels are commonly specified, over

a wide range of weld energy inputs and weld type.

(c) Gaining insights, where possi ble, into the factors

affecting response to welding and the prop erti es a chie ved.

Three steels were used in the investigation; BS4360:50D and Euronorm 25-72

Fe510 DD, both of which were normalised gr ades , and RQT 500, a BSC proprietar y

quenc hed and tempered ste el, of appro ximat ely the same compo sitio n as

BS4 360 :50D . Welds were carried out at heat inputs of 2 and 5 kJ/m m, using

single V, double V and K prep arat ions , together with electroslag wel ds at

30-50 kJ/mm , dependi ng on plate thi ckn ess . The majori ty of the work was

carried out on 40 mm thick wel ds , with some compa rativ e tests at 25 and 60 mm.

The small scale tests used (Charpy V, 10 χ 10 mm COD, full thi ckness C O D ) , all

ranked the steels in the same ord er. The BS436 0:50D pla te , at 40 mm thick,

performed bes t, with RQT 500 second and Fe510DD third. The major factor

affecting HAZ properties, at least in the sample plates tested, appeared to be

the carbon equivalent value

(CEV),

which was about 0.4% in the 50D and RQT 500

plates, and 0.5% in the Fe510D D samp le. The Charpy V and COD tests showed

that,

in the coarse grained region of the HAZ , occasi onal low results could be

obtained, but, in wide plate tests at -30 and -40°C, in the presence of

9 χ 90 mm fatigued surface not che s, located in the HAZ , stresse s of the order

of plate yield lev el, and overall strai ns ranging from 1.1 to 7.5, depen ding on

the ste el, were sustained before fr actu re. Two of the plate results were

analysed, by the procedures given in PD

6493,

and it was shown tha t, as in

previous experience with weld metals, tolerable defect predictions in HAZ's

have an in-built safety fact or, in these two cases , of 2-3.

With respect to the methodology of testing, the correct placement of HAZ

notches presented the only significant problems . There was very evident

scattter in the results from all tests, although this was also a feature of

tests on some of the parent plates.

Sub-size COD tes ts, 10 mm square in cross sec tio n, giving more accurat e

indications of fracture initiation resistance, and the generation of

microstructures typical of

HAZ's,

by simulation of thermal cycles derived from

the test we ld s, were shown to have potent ial in research wo rk, although of

limited direct applicability to practical si tuatio ns.

A large body of thermal cycle data was generated from thermocouples embedded in

selected weld

HAZ's.

Good agree ment with publish ed work was obt ain ed, and

suggestions were made for modifying some of the physical constants used in

theoretical predicitions.


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FR 62-10 822 7210.KA/804

CONTENTS PAGE

1. INTRODUCTION 1

1.1 The Nature of the Heat Affected Zone 1

1.2 Importance of Heat Affected Zone Properties 3

2.

EXPERIMENTAL PROGRAMME 8

2.1 Aims and Objectives of the Work 8

2.2 Steels Used in the Test Programme 8

2.3 Weld Procedures 10

2.4 Test Methods 11

3. TEST RESULTS 31

3.1 Charpy V Impact Test Results 31

3.2 Sub-size (10 χ 10 mm Section) COD Test Results 33

3.3 Full Thickness COD Tests 33

3.4 Wide Plate Test Results 35

3.5 Test Results from Simulated HAZ's 37

4.

GENERAL DISCUSSION 75

4.1 Appraisal of Test Methods 75

4.2 Relevance of Simulative Studies 77

4.3 Effect of Steel Grade on Results Obtained 77

5. CONCLUSIONS 80

6. REFERENCES 81

TABLES

FIGURES

APPENDIX

111


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FR 62-10 822 7210.KA/804

LIST OF TABLES

2.1 Chemical Analyses of Plates Tested

2.2 Mechanical Properties of Plates Tested

2.3 Sub-Division of Weld Test Plate

3.1 Matrix of Tests Carried Out

3.2 HAZ Charpy V Test Criteria - BS4360:50D (40 mm Thick)

3.3 HAZ Charpy V Test Criteria - BS4360:50D (25 and 60 mm Thick)

3.4 HAZ Charpy V Test Criteria - Euronorm 25-72 Fe510DD (40 mm Thick)

3.5 HAZ Charpy V Test Criteria - RQT 500 (40 mm Thick)

3.6 HAZ Charpy V Test Criteria - RQT 500 (25 mm Thick)

3.7 Proportions of Weld Metal/HAZ/Plate Adjacent to Notch of Subsurface

Single and Double V Weld Charpy Tests (Plate/HAZ Boundary Taken as

Visible Aci)

3.8 0.1 mm COD Transition Temperatures (10 mm Square Transverse

40 mm Welds)

3.9 Summary of COD Test Results

3.10 Wide Plate Test Results

3.11 Wide Plate Test - Notch Locations and Dimensions


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FR 62-10 822 7210. KA/80 4

LIST OF FIGURES

1.1 Heat Affecte d Zone Thermal Cycl es - 5 kJ/mm

1.2 Weld Macro graph (5 kJ/mm , Single V) Showing Reheated Heat

Affected Zones

2.1 Microstr uctures of As-received Normalised Plates

(a) Fe51 0DD, 40 mm Thick

(b) BS4 360 :50 D, 25 mm Thick

(c) BS43 60:5 0D, 40 mm Thick

(d) BS4 360 :50 D, 60 mm Thick

2.2 Microstr uctures of As-received Quenched and Tempered Plate s,

RQT 500, 25 mm Thick

2.3 Typi cal 5 kJ/mm K Weld Proce dure

2.4 Typic al 2 kJ/mm Double V Weldin g Proce dure

2.5 Typical Electroslag Weld Procedure

2.6 Typic al 25 mm Single Bevel Weld Proce dure

2.7 Wid e Plate Test Format - Showing Locati ons of Instr ument ation

3.1 41 J Impact Transition Temperatu res - 2 kJ/mm Welds

3.2 41 J Impact Trans ition Tem pera tur es - 5 kJ/mm Welds

3.3 41 J Impact Trans ition Tem pera tur es - 2 and 5 kJ/mm K We ld s,

25 and 60 mm Thick

3.4 Shift in 41 J Impact Tran sitio n Temp erat ures - 2 kJ/mm Wel ds

3.5 Shift in 41 J Impact Tran siti on Tem pera tur es - 5 kJ/mm Wel ds

3.6 Shift in 41 J Impact Tran siti on Temp erat ure - 2 and 5 kJ/mm

K Welds, 25 and 60 mm Thick

3.7 Parent Pla te, Charpy V Impact Energy Curve Showing Scatter in

Results

3.8 HAZ Charpy V Impact Energy Curves Showing Scatter in Resul ts

3.9 Specimen Locat ions and Notch Posi tion s Charpy and 10 χ 10

COD Specimens

3.10 10 mm Square COD Result s Showing Scatter

3.11 Parent Plate Full Thick ness COD Tests Tra nsv ers e, 40 mm Thick

3.12 Comp aris on of COD Result s - 40 mm Plate s

3.13 COD Res ult s, 25 and 60 mm BS4360 :50D Plates

3.14 COD Res ult s, 25 mm RQT 500 Plates

3.15 COD Result s on Weld D4W2, 50D, Double V, 2 kJ/mm

3.16 Full Thick ness COD Results - Fe51 0DD, 5 kJ/mm Κ Weld

3.17 Full Thickn ess COD Results - BS4 360: 50D , 5 kJ/mm Κ Weld

3.18 Full Thickn ess COD Results - Fe5 10D D, Electr oslag Weld (_50 kJ/mm)

v u


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FR 62-10 822 7210.KA/804

3.19(a) Effect of Peak Temperature in Simulation 5 kJ/mm Cycle -

BS4360:50D

Effect of Tempering Cycle - BS4360:50D

Effect of Peak Temperature in Simulation 5 kJ/mm Cycle -

Fe510DD

Effect of Tempering Cycle - Fe510DD

Effect of Peak Temperature in Simulation 5 kJ/mm Cycle - RQT 500

Effect of Tempering Cycle - RQT 500

Effect of Peak Temperature - 2 kJ/mm Cycle - BS4360:50D

Effect of Peak Temperature - 2 kJ/mm Cycle - Fe510DD

Effect of Peak Temperature - 2 kJ/mm Cycle - RQT 500

Simulated HAZ Microstructures - BS4360:50D

(At

8

_

5

= 50 s,

approximating to a 5 kJ/mm heat input)

3.26 Simulated HAZ Microstructures - Fe510DD (Atg-s = 50 s, approximating

to a 5 kJ/mm heat input)

3.27 Simulated HAZ Microstructures - RQT 500 (At

8

_

5

= 50 s, approximating


3.28 Simulated HAZ Microstructures - BS4360:50D (At

8

_

5

= 20 s,

approximating to a 2 kJ/mm heat input)

3.29 Simulated HAZ Microstructures - Fe510DD (At

8

-s = 20 s, approximating


3.30 Simulated HAZ Microstructures - RQT 500

(At

8

_

5

= 20 s, approximating


4.1 Comparison of Charpy V and 10 mm Square COD Transition Temperatures

3 .

3 .

3 .

3 .

3 .

3 .

3 .

3 .

3 .

1 9 b )

2 0 a )

, 2 0 b )

2 1 a )

2 1 b )

2 2

2 3

24

, 2 5

V i l i


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MESURE DE LA TENACITE DANS LA ZONE DE TRANSITION

DES ACIERS DE CONSTRUCTION SOUDES

Accord C.E.C.A. No. 7210.ΚΑ/804

SOMMAIRE

Etat donné l'importance croissante des propriétés de la zone de transition dans les

cahiers des charges ainsi que pour le comportement en service des aciers utilisés en

construction et soumis à des conditions sévères, une étude a été menée dans le but de:-

(a) examiner de manière critique les méthodes dont on dispose à

l heure actuelle afin de déterminer la ténacité à la rupture

des zones de transition, en particulier lorsqu'il s agit

d'établir l'aptitude à l'emploi et de pouvoir prédire la pe r-

formance éventuelle en service des zones de transition, à

partir d'essais conduits sur une échelle limitée, par compa-

raison avec les essais pratiqués sur des plaques larges

présentant des zones de transition en entailles.

(b) établir les propriétés qu'il est possible d'obtenir dans les

zones de transition de certains aciers de construction

utilisés dans des cas où il est courant qu'on exige pour les

zones de transition des caractéristiques p oussée s, et ce pour

un large éventail d'apports d'énergie et de types de soudures.

(c) parvenir dans la mesure du possible à une meilleure compréhen-

sion des facteurs qui influent sur la réponse au soudage, ainsi

que des propriétés obtenues.

Trois aciers ont été utilisés pour cette étude, à savoir BS4360:50D et Euronorm 25-72

Fe510DD,deux nuances normal isées, et RQT 500 , un acier propre à la BSC, qui a subi trempe

et revenu et dont la composition est quasi identique à celle de BS4360:50D. Des soudures

ont été exécutées pour des apports de chaleur de 2 et 5 kJ/mm, pour des préparations de

joints chanfreinés en V, en X et en K, ainsi que des soudures sous laitier à 30-50 kJ/mm

selon l'épaisseur de la plaque. La majeure partie de l étude a porté sur des soudures

de 40 mm d'épaisseur, et certains essais comparatifs ont été pratiqués à 25 et à 60 mm.

Les essais sur une petite échelle (Charpy V, déplacement de l'ouverture de la fissure

10 xlO mm, déplacement sur épaisseur complète) ont tous donné un même ordre pour les

aciers, le meilleur étant la plaque de 40mm d'épaisseur de BS4360:50D, suivie de

RQT 500, Fe510DD arrivant en troisième lieu. En ce qui concerne les propriétés de la

zone de transition, du moins pour les échantillons soumis aux essais, il semble que le

facteur le plus important soit l'équivalent en carbone, qui était de l ordre de 0,4% pour

les plaques en 50D et RQT 500 , et de 0,5% pour l'échantillon de Fe510DD. Les essais

Charpy V et essais de déplacement de l'ouverture des fissures ont révélé que des résultats

faibles pouvaient parfois être obtenus dans la région à gros grain de la zone de tran si-

tion;

toutefois lors des essais sur plaques larges pratiqués à -30 et -40°C, en présence,

dans la zone de transition, d'entailles en surface de 9 χ 90mm soumises à la fatigue, on

a enregistré avant la rupture des sollicitations de

l ordre

de celles se produisant à

la limite élastique, et des taux de travail allant globalement de 1,1 à 7,5 selon la

nature de l'acier. Deux séries de résultats obtenus sur ces plaques ont été analysées

selon les méthodes indiquées en PD 6493, et ces analyses ont montré que comme on le savait

déjà pour les métaux d'apport, les prévisions de défauts tolerables dans les zones de

transition comportent un facteur de sécurité qui, pour ces deux cas, était de 2 - 3.


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En ce qui concerne la méthodologie des essais le seul problème appréciable a été de

trouver l'endroit approprié pour les entailles des zones de transition. On a enregist:

pour les résultats de tous les essais une très nette dispersion, qui avait d'ailleurs

déjà remarquée lors des essais sur certaines plaques de base.

On a pu voir que les essais de déplacement de l'ouverture des fissures sur de très pet:

échantillons ayant une section transversale de 10 mm de côté, qui donnent des indicate

plus précises sur la résistance à l'amorce de rupture, ainsi que la simulation des cyc

thermiques dérivés des soudures d'essai, qui engendre des microstructures caractéristii

des zones de transition, peuvent donner des résultats intéressants pour la recherche; ;

contre les possibilités d'application directe dans des cas pratiques sont limitées.

Des thermocouples pénétrant dans des zones de transition sélectionnées ont fourni un imi

tant volume d'informations sur les cycles thermiques. Les résultats ont bien concordé

avec les travaux déjà publiés et il a été possible de suggérer des modifications des

constantes physiques utilisées pour les prédictions théoriques.


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FR 62-10 822

SOMMAIRE

1, INTRODUCTION

1.1 Nature de la zone de transition

1.2 Importance des propriétés de la zone de transition

7210.KA/804

PAGE

1

1

3

8

8

8

10

11

2.

PROGRAMME EXPERIMENTAL

2.1 Buts et objectifs des travaux

2.2 Aciers utilisés pour le programme d essais

2.3 Modes de soudage

2.4 Méthodes d essai

3. RESULTATS DES ESSAIS 3J

3.1 Résultats des essais d impact Charpy V 31

3.2 Résultats des essais de déplacement de l ouverture des

fissures sur échantillons de section 10 χ 10 mm 33

3.3 Essais de déplacement de l ouverture des fissures à pleine

épaisseur 33

3.4 Résultats des essais sur plaques larges 35

3.5 Résultats des essais pratiqués sur des zones de transition

simulées 37

4. DISCUSSION GENERALE 75

4.1 Evaluation des méthodes d essai 75

4.2 Intérêt d études simulatives 77

4.3 Effet de la nuance d acier sur les résultats obtenus 77

5. CONCLUSIONS 80

6. REFERENCES 81

TABLES

CHIFFRES

ANNEXE

xi


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LISTE DES TABLEAUX

2.1 Analyses chimiques des plaques soumises aux essais

2.2 Propriétés mécaniques des plaques soumises aux essais

2.3 Sous-division des plaques soumises aux essais de soudure

3.1 Matrice des essais pratiqués

3.2 Critères essais Charpy V sur zone de transition - BS4360:50D (ép. 40mm)

3.3 Critères essais Charpy V sur zone de transition - BS4360:50D (ép. 25 et 60 mm)

3.4 Critères essais Charpy V sur zone de transition - Euronorm 25-72 Fe510DD (ép.

40 mm)

3.5 Critères essais Charpy V sur zone de transition - RQT 500 (ép. 40 mm)

3.6 Critères essais Charpy V sur zone de transition - RQT 500 (ép. 25 mm)

3 η Proportions de métal d'apport/zone de transition/plaque au voisinage de

l'entaille, essais Charpy sur soudures en V et en X, en-dessous de la sur-

face (limite plaque/zone de transition supposée être située à la transformation

visible).

3.8 Températures de transition, déplacement de l'ouverture des fissures 0,1 mm

(soudures transversales 40 mm, 10mm de côté)

3.9 Sommaires des résultats des essais de déplacement de l'ouverture des fissures

3.10 Résultats des essais sur plaque large

3.11 Essai sur plaque large - emplacement et dimensions des entailles

Xlll


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LISTE DES FIGURES

1.1 Cycles thermiques, zone de transition - 5 kJ/mm

1.2 Macrographie de soudure (5 kJ/mm , en V) montrant les zones de

transition réchauffées

2.1 Microstructures des plaques brutes de normalisation

(a) Fe510DD, ép. 40 mm

(b) BS4360:50D, ép. 25 mm

(c) BS4360:50D, ép. 40 mm

(d) BS4360:50D, ép. 60 mm

2.2 Microstructures des plaques brutes de trempe et revenu,

RQT 50 0, ép. 25 mm

2.3 Mode opératoire typique 5 kJ/mm, soudure en K

2.4 Mode opératoire typiq ue, 2 kJ/mm, soudure en X

2.5 Mode opératoire typique, soudure sous laitier

2.6 Mode opératoire typiq ue, soudure un chanfrein, 25 mm

2.7 Emplacement des instruments pour les essais sur plaque large

3.1 Températures de transition, impact 41 J - soudures 2 kJ/mm

3.2 Températures de transition, impact 41 J - soudures 5 kJ/mm

3.3 Températures de transition, impact 41 J - soudures de 2 et 5 kJ/mm , en K,

ép. 25 et 60 mm

3.4 Décalage des températures de transition, impact 41 J - soudures 2 kJ/mm

3.5 Décalage des températures de transition, impact 41 J - soudures 5 kJ/mm

3.6 Décalage des températures de transition, impact 41 J - soudures 2 et 5 kJ/mm

en K, ép. 25 et 60 mm

3.7 Plaque de bas e, courbe Impact Charpy V - Energie montrant la dispersion des

résultats

3.8 Courbes Impact Charpy V - Energie pour les zones de transition, montrant

la dispersion des résultats

3.9 Emplacement des êprouvettes et des entail les, éprouvetttes Charpy et

de déplacement d'ouverture des fissures 10 χ 10

3.10 Résultats déplacement de l'ouverture des fissures, 10mm de cô té, montrant

la dispersion

3.11 Essais transversaux déplacement de l'ouverture des fissures pleine épaisseur

de la plaque de ba se, ép. 40 mm

3.12 Comparaison des résultats de déplacement de l'ouverture des fissures - plaques

de 40 mm

3.13 Résultats déplacement de l'ouverture des fissures, plaques BS4360:50D, 25 et

60 mm

3.14 Résultats déplacement de l'ouverture des fissures, plaques RQT 50 0, 25 mm

3.15 Résultats déplacement de l'ouverture des fissures sur soudure D4W2, 50D,

en X, 2 kJ/mm

3.16 Résultats déplacement de l'ouverture des fissures pleine épaisseur - Fe510DD,

5 kJ/mm. soudure en K

3.17 Résultats déplacement de l'ouverture des fissures pleine épaisseur - BS4360:50D,

5 kJ/mm, soudure en K

3.18 Résultats déplacement de l'ouverture des fissures pleine épaisseur - Fe510DD,

soudure sous laitier (- . 50 kJ/mm)


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FR 62-10 822 7210.KA/804

3.19(a) Effet de la température de pointe, simulation cycle de. 5 kJ/mm -

BS4360:50D

3.19(b) Effet du cycle de revenu - BS4360:50D

3.20(a)

Effet de la température de pointe, simulation cycle de 5 kJ/mm -

FE510DD

3.20(b) Effet du cycle de revenu - Fe5lODD

3.21(a) Effet de la température de pointa, simulation cycle de 5 kJ/mm - RQT 500

3.21(b) Effet du cycle de revenu - RQT 500

3.22 Effet de la température de pointe - cycle 2 kJ/mm - BS4360:50D

3.23 Effet de la température de pointe - cycle 2 kJ/mm - Fe510DD

3.24 Effet de la température de pointe - cycle 2 kJ/mm - RQT 500

3.25 Microstructures zones de transition simulées - BS4360:50D (kt„ = 50 s,

soit apport de chaleur d'environ 5 kJ/mm)

3.26 Microstructures zones de transition simulées - Fe510DD (i t „ , = 50 s,


3.27 Microstructures zones de transition simulées - RQT 500 (&t„ = 50 s,


3.28 Microstructures zones de transition simulées - BS4360:50D (ût„_

c

. = 20 s,


3.29 Microstructures zones de transition simulées - Fe510DD (ût„ _ = 20 s,


3.30 Microstructures zones de transition simulées - RQT 500 (At

fi

_,. = 20 s,


4.1 Comparaison des températures de transition Charpy V et déplacement de

l'ouverture des fissures 10mm de côté

X V I


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Messung der Festigkeit in der wärmebeeinflußten Zone der

geschweißten Konstruktionsstähle

EKSG Vertrag Nr. 7210.KA/804

Zusammenfassung

In Anbetracht auf die zunehmende Bedeutung der Eigenschaften der wärme-

beeinflußten Zone (HAZ), die in den Spezifikationen vorgeschrieben werden,

und den hohen Anforderungen, die an das Einsatzverhalten der Stähle in

Konstruktionsverwendungszwecken gestellt werden, wurde eine Untersuchung

durchgeführt, die die folgenden Ziele hatte:

a. Kritische Prüfung der zur Zeit einschlägigen Verfahren für die Unter-

suchung der Bruchfestigkeit der HAZ und zwar unter besonderer Bezugs-

nahme auf die Bewertung der Eignung für den Zweck und auf die Fähigkeit,

die potentielle Leistung der HAZ im Einsatz mit den Ergebnissen der

Prüfungen im kleinen Umfang im Vergleich zu den breiten Metallplatten-

prüfungen vorherzusagen, die HAZ Kerben enthielten.

b. Bewertung der Eigenschaften, die in den HAZ von einigender Baustähle

erreicht werden, denn für diese zur Zeit eingesetzten Stähle wird

häufig ein hohes Eigenschaftsniveau vorgeschrieben. Die Bewertung

wurde für viele verschiedene Schweißenergiezufuhren und Schweißtypen

durchge

führt.

c. Wo möglich, Gewinnung von Einblicken in die Faktoren, die die Reaktion

auf die Schweißung und die gewonnenen Eigenschaften beeinflußen.

In der Untersuchung wurden drei Stähle eingesetzt: der BS4360:50D und der

Euronorm 52-72 Fe5lODD, die beide eine normalisierte Güte hatten, und der

RQT 500, der ein vergüteter Markenstahl der British Steel Corporation ist

und ungefähr die gleiche Zusammensetzung wie der BS4360:50D hat. Die

Schweißungen wurden bei einer Wärmezufuhr von 2 und 5 kJ/mm unter Einsatz

von einfachen V, doppelten V und K Formen zusammen mit Elektroschlacken-

schweißungen von 30 - 50 kJ/mm durchgeführt, das hing von der Metallplatten-

stärke ab. Der größte Teil der Forschung wurde auf 40 mm starken Schweis-

sungen durchgeführt, aber man machte auch einige vergleichende Prüfungen auf

25 und 60 mm starken Platten.

Man konnte mit den im kleinen Umfang gemachten Prüfungen (Charpy V, ΙΟ χ 10 mm

COD,

volle Stärke COD) alle Stähle in der gleichen Anordnung einstufen. Die

BS4360:50D Metallplatte mit einer Stärke von 40 mm hatte die beste Leistung,

die RQT 500 kam an zweiter Stelle und die Fe5lODD an dritter. Es schien,

alsob der bedeutendste Faktor, der die HAZ Eigenschaften zumindest in den

geprüften Probenmetallplatten beeinflußte, der Kohlenstoffäquivalentwert

(C.EV) war, der in den 50D und den RQT 500 Platten bei ungefähr 0,4% und

in der Fe510DD Probe bei 0,5% lag. Die Charpy und COD Prüfungen zeigten,

daß in dem grob gekörnten Bereich der HAZ zuweilen niedrige Ergebnisse

gewonnen werden konnten, aber in den breiten Metallplattenprüfungen bei

-30 und -40° C, wo ermüdete 9 χ 90 mm Oberflächenkerben in der HAZ vorlagen,

wurden Beanspruchungen in der Anordnung des Plattennachgebeausmaßes und

der Gesamtverzerrungen vor dem Bruch bestätigt, die von 1,1 bis zu 7,5

reichten, was von dem Stahl abhängig war. Zwei der Metallplattenergebnisse

wurden gemäß dem im PD 6493 angegebenen Verfahren analysiert, und man konnte

xvii


24/204

FR 62-10 822 7210.KA/804

mit den aus früheren mit geschweißten Metallen gewonnenen Erfahrungen zeigen,

daß zulässige Defektvorhersagen in den HAZ einen eingebauten Sicherheits

faktor haben, der in diesen beiden Fällen bei 2 - 3 lag.

Hinsichtlich der Prüfmethodologie war nur die richtige Anordnung der HAZ

Kerben das einzige bedeutende Problem. Man gewann eine sehr deutliche

Streuung in den Ergebnissen in allen Prüfungen, aber dies war auch eine

charakteristische Eigenschaft der Prüfungen in einigen der Ausgangsmetall

platten.

Untergrößen COD Prüfungen wurden auf Proben mit einem Querschnitt von 10 mm

2

gemacht, um genauere Hinweise auf den Bruchanfangswiderstand und die für die

HAZ typischen Mikrogefüge zu gewinnen. Prüfungen wurden durch Simulierung

der thermischen Zyklen gemacht, die aus den Prüfschweißungen abgeleitet

worden waren, und man konnte damit zeigen, daß sie ein Potential in der

Forschungsarbeit haben, obwohl sie nur begrenzt direkt in praktischen

Situationen anwendbar sind.

Ein großer Teil der thermischen Zyklusdaten wurde durch die in ausgewählten,

geschweißten HAZ eingekapselten Thermoelementen erzeugt. Man gewann gute

Übereinstimmung mit den veröffentlichten Forschungsarbeiten, und Vorschläge

wurden für die Modifizierung einiger der physischen Konstanten gemacht,

die in theoretischen Vorhersagen benutzt werden.

x v m


25/204


26/204


27/204

FR 62-10 822 7210.KA/804

Aufstellung der Tabellen

2.1 Chemische Analyse der geprüften Metallplatten

2.2 Mechanische Eigenschaften der geprüften Metallplatten

2.3 Aufgliederung der geschweißten Prüfmetallplatten

3.1 Matrix der durchgeführten Prüfungen

3.2 HAZ Charpy gegen die Prüfkriterien - BS4360:50D (40 mm stark)

3.3 HAZ Charpy gegen die Prüfkriterien - BS4360:50D

(25 und 60 mm stark)

3.4 HAZ Charpy gegen die Prüfkriterien - Euronorm 25-72 Fe5lODD

(40 mm stark)

3.5 HAZ Charpy gegen die Prüfkriterien - RQT 500 (40 mm stark)

3.6 HAZ Charpy gegen die Prüfkriterien - RQT 500 (25 mm stark)

3.7 Verhältnis des Schweißmetalls:HAZ:Metallplatte neben den

unter der Oberfläche befindlichen Kerben, einfache und doppelte

V geschweißte Charpy Prüfungen (Metallplatten/HAZ-Grenze von dem

sehbaren Ac^ genommen)

3.8 0,1 mm COD Übergangstemperaturen (IO mm

2

diagnonale 40 mm

Schweißungen)

3.9 Zusammenfassung der COD Prüfergebnisse

3.10 Prüfergebnisse der breiten Metallplatten

3.11 Prüfung der breiter Metallplatten - Kerbstellen und Dimensionen

xxi


28/204


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FR 62-10 822 7210.KA/804

Aufstellung der Abbildungen

1.1 Thermische Zyklen der wärmebeeinflußten Zone - 5 kJ/mm

1.2 Geschweißtes Makrodiagramm (5 kJ/mm, einfaches V ) ,

gezeigt werden die wieder erwärmten, wärmebeeinflußten

Zonen

2.1 Mikrogefüge der wie erhaltenen, normalisierten

(a) Metallplatten, Fe510DD, 40 mm stark

(b) BS4360:50D, 25 mm stark

(c) BS4360:50D, 40 mm stark

(d) BS4360-.5CD, 60 mm stark

2.2 Mikrogefüge der wie erhaltenen, vergüteten

Metallplatten, RQT 50O, 25 mm stark

2.3 Typisches 5 kJ/mm Schweißverfahren

2.4 Typisches 2 kJ/mm, doppeltes V Schweißverfahren

2.5 Typisches Elektroschlackenschweißverfahren

2.6 Typisches 25 mm einfaches Schrägschweißverfahren

2.7.

Breites Metallplattenprüfverfahren, gezeigt wird die

Anordnung der Instrumentierung

3.1 41 J Aufschlagsübergangstemperaturen - 2 kJ/mm Schweißungen

3.2 41 J Aufschlagsübergangstemperaturen - 5 kJ/mm Schweißungen

3.3 41 J Aufschlagsübergangstemperaturen - 2 und 5 kJ/mm K Schweißungen,

25 und 60 mm stark

3.4 Verschiebung in den 41 J Aufschlagsübergangstemperaturen -

2 kJ/mm Schweißungen

3.5 Verschiebung in den 41 J Aufschlagsübergangstemperaturen -

5 kJ/mm Schweißungen

3.6 Verschiebung in der 41 J Aufschlagsübergangstemperatur -

2 und 5 kJ/mm K Schweißungen, 25 und 60 mm stark

3.7 Ausgangsmetallplatte, Charpy V Aufschlagsenergiekurve,

gezeigt wird die Streuung in den Ergebnissen

3.8 HAZ Charpy V Aufschlagsenergiekurven, gezeigt wird die Streuung

in den Ergebnissen

3.9 Probeanordnungen und Kerbstellen, Charpy und 10 χ 10 COD Proben

3.10 10 mm

2

COD Ergebnisse, gezeigt wird die Streuung

3.11 Ausgangsmetallplatte, volle Stärke, diagonale COD Prüfungen, 40 mm stark

3.12 Vergleich der COD Ergebnisse - 40 mm Metallplatten

3.13 COD Ergebnisse, 25 und 60 mm, BS4360:50D Metallplatten

3.14 COD Ergebn isse, 25 mm, RQT 500 Metallplatten

3.15 COD Ergebnisse der Schweißung

D4W2,

50D, doppeltes V, 2 kJ/mm

3.16 COD Ergebnisse der vollen Stärke - Fe5lODD, 5 kJ/mm Κ Schweißung

3.17 COD Ergebnisse der vollen Stärke - BS4360:50D, 5 kJ/mm Κ Schweißung

3.18 COD Ergebnisse der vollen Stärke - Fe510DD, Elektroschlacken-

schweißung (̂ 50 kJ/mm)

xxiii


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FR 62-10 822 7210.KA/804

Aufstellung der Abbildungen (Forts.)

3.19 Auswirkung der Spitzentemperatur in der Simulierung, 5 kJ/mm

(a) Zyklus - BS4360:50D

3.19 Auswirkung des Vergütungszykluses - BS4360:50D

(b)

3.20 Auswirkung der Spitzentemperatur in der Simulierung, 5 kJ/mm Zyklus

(a) Fe510DD

3.20 Auswirkung des Vergütungszykluses - Fe510DD

(b)

3.21 Auswirkung der Spitzentempratur in der Simulierung, 5 kJ/mm

(a) Zyklus - RQT 500

3.21 Auswirkung des Vergütungszykluses - RQT 500

(b)

3.22 Auswirkung der Spitzentemperatur - 2 kJ/mm Zyklus - BS4360:50D

3.23 Auswirkung der Spitzentemperatur - 2 kJ/mm Zyklus - Fe510DD

3.24 Auswirkung der Spitzentemperatur - 2 kJ/mm Zyklus - RQT 500

3.25 Simulierte HAZ Mikrogefüge - BS4360:50D, (Ate~5

= 5 0 s

'

a n

eine

eine 5 kJ/mm Wärmezufuhr angeglichen)

3.26 Simulierte HAZ Mikrogefüge - Fe5l0DD (Atg~5 = 50 s, an eine

5 kJ/mm Wärmezufuhr angeglichen)

3.27 Simulierte HAZ Mikrogefüge - RQT 500 (Ate~5 = 50 s, an eine


3.28 Simulierte HAZ Mikrogefüge - BS4360:50D,

(àtg-

5

=

20 s, an

eine 2 kJ/mm Wärmezufuhr angeçlichen)

3.29 Simulierte HAZ Mikrocefürge - Fe510DD (At8~5 = 20 s, an

eine 2 kJ/mm Wärmezufuhr angeglichen)

3.30 Simulierte HAZ Mikrogefüge - RQT 500 (Ate~5 = 20 s, an eine


4.1 Vergleich zwischen den Charpy V und 10 mm

2

C0D Übergangs

temperaturen

xxiv


31/204

FR 62-10 822 7210. KA/80 4


MEASUREMENT OF TOUGHNESS IN THE HEAT AFFECTED ZONE

OF WELDED STRUCTUR AL STEELS

ECSC Agreement No. 7210.KA/804

FINAL TECHNICAL REPORT

1. INTRODUCTION

1.1 The Natu re of the Heat Affec ted Zone

The visible weld heat affected zone (HAZ) is by definition, that part of the

weld joint lying between the fusion boundary with the weld metal and that part

of the structure retaining, at the optical microscopy level, the original

as-delivered micro stru cture . Within this narrow band, a few millimet res wi de ,

the steel will have been subjected to one or more thermal cycles due to heating

by the individu al weld beads laid do wn. The nature of these thermal cycles at

a given location in the HAZ is a function of the position relative to the weld

bead, and the factors affecting he at flow (thickn ess, initial steel

temperature, weld energy input). The family of curves determ ined in this

programme are shown in Fig. 1.1. In a multirun weld, metal at a given point

may be subjected to a significant number of consecutive cycles with differing

peak tempe rature s and cooling ra tes . The reheated areas of HAZ's can often be

observed in macrogra phs of multirun we lds , Fig. 1.2.

Within this narrow band of material, there exists a continuum of heating

cycles,

decre asing in peak tem perat ure and overall cooling rate , from the

fusion boundary to the outer HAZ edge , giving rise to micr ost ruct ures which

change continuously over this region. A number of general zones, within the

visible HAZ, can be identified, the extent and nature of which will vary with

steel compositi on, the thermal cycle sustained and, to a variable e xtent, the

original microstructure:-

(a) Grain Coars ened HAZ

In this region, adjacent to the fusion boundary, the original microstructure

has been reaustenitised at temperatures and retention times sufficient for

marked grain growth to occ ur. The aust enit ic grain size attained will vary

according to a number of factors :-

Distance from the fusion boundary, being greatest

immediately adjacent to it.

Weld energy input, as a function of peak temperature

and relative retention time.

Original grain size; under the non-equilibrium

conditions,

finer austenite grain sizes may result

from finer original microstructures.

Steel composition, particularly with respect to the

presence of refractory carbides or nitrides which

are difficult to dissolve in the available timescale

at high temperature, thereby exercising a pinning

effect on the austenite grain bounda ries. Compositi on

may also affect the austenitising temperature itself

and, hence, the extent of coarsening.

The subsequent transformation products will be defined by the cooling rate,

over the transforma tion ran ge, and the composition and grain size of the prior

austenite, and so may differ in detail from steel to steel, depending on the

prior thermal histor y. In gene ral, howeve r, the grain coarsened region

will consist of a variable mixture of constituents

1

·

1

. These will include one

or more of the following, in order of decreasing transformation temperature:

proeutectoid ferrite at prior austenite grain boundaries or in transgranular


32/204

FR 62-10 822 7210 .KA/8 04

Widmanstatten plate form; areas of high carbon content ranging from pearlite to

morpho logies in which carbides precipit ate as rods or spher oids, bainitic

colon ies in which the ferrite pla tes may interlock or grow side-b y-sid e in a

manner resembling classica l upper baini te, with plate widths dependent on

transform ation temper ature; lower bainite; marte nsit e.

(b) Grain Refined Regio n

As the peak temperatures fall, the situation will eventually arise where the

HAZ has been reheated just into the auste nitis ing tempe ratur e ra nge , bearing in

mind the rapid heating rate and short retention time . This will give rise to

very fine aust eni te, which coupled with the slower cooling rates at the

increased distance from the fusion boundary, transforms to a fine equiaxed

ferrite matrix with small areas of ferrite/carbide aggregate mixtures.

(c) Intercritical Region

In this portion of the heat affected zone, the Ac2 is only exceeded in the high

carbon constituents of the original microstructure, and the original ferrite is

largely untouched . Some dilution of the high carbon austenite regions may

occur by diss olut ion of the surrou nding fer rit e, and the subse quent transforme d

stru ctur e will depend on their compo sitio n and the local cooling rate. As

the peak temperature progressively decreases, the size of the reaustenitised

region s dec rea ses , and the limit of the visible HAZ is defined by a region in

which the structure of the pre-e xisti ng p earli te is merely degraded by some

extent of spheroidisation.

(d) Subcriticai Region

Out side the visibl e HAZ is a relativ ely broad zone in which the temp eratur e

peaks are low, but the heating time s are long . This region was of great

importance with C and CMn steels, since it was possible for nitrogen strain

ageing to occur, and was, indeed, the reason for the development of the

original Wells wide plate test, to investigate the effect of the consequent

loss in cleavage fracture resistance, in full thickness welds.

The two principal factors needing to be controlled have been shown to

b e

( 1 . 2 Λ . 3 ) . :

The ferrite grain size of the parent material.

The free or interstitial nitrogen content.

These factors are adequately dealt with in the more modern fine grained CMnNb

and CMnNbAl steels, and nitrogen strain ageing is no longer the potential

problem it was.

The various regions described above merge into each other and it would be

impossible in most instances to decide on a distinct boundary between them.

In practice, there is no great necessity to do so, since, so far as fracture

resistance is concerned, it is the grain coarsened region which show the

greatest degree of degradation from the original properties.

The extreme heterogeneity of even the HAZ adjacent to a single run weld, to

which must be added the variability of tempering by subsequent runs, in a

multirun situation, delineates the problem facing the research worker

investigating the properties, or even the microscopy, of real welds. The

process zone associated with the smallest mechanical test piece is bound to

encompass regions of markedly different microstructures and properties, tending

to blur any relationships sought between structure and properties and,

possibly, to introduce scatter in repeat tests taken from nominally the same

position with reference to the weld.

To some degree, the original steel is itself heterogeneous with variable

structures from surface to centre and centre to edge in plate products, and

between regions of differing thickness in rolled sections. Segregation of

various alloying elements occurs, and its effects on transformation can be

clearly seen in many HAZ macrographs. Non-metallic inclusions are always

directionally aligned, to some extent, and will ensure that, in particular,


33/204

FR 62-10 822 7210.KA/804

fracture toughness may show marked anisotropy, between longitudinal, transverse

and short transverse specimen orientations. Even when a single orientation is

taken, many types of test result show a wide normal distribution. It would be

unwise to expect a similar population of HAZ results to exhibit more convenient

characteristics.

The foregoing, therefore, presents some of the basic difficulties involved in

any attempts to characterise HAZ properties or to compare the HAZ performance

of different steels. The main aim of this programme was to assess the extent

of these problems, in real welds, when carrying out the types of fracture

toughness tests employed by the fabrication industry. This would provide

information enabling a judgement to be made on how much significance should be

attached to very limited programmes of testing, often carried out by

fabricators or even classification societies under conditions less controlled

than those of the research laboratory.

1.2 Importance of Heat Affected Zone Properties

The overall aim in a welded joint, with respect to fitness-for-purpose, must be

to obtain a joint in which the three component parts - parent steel, HAZ and

weld metal, are each capable of withstanding the demands which the application

places upon them, bearing in mind the probable size and location of defects

either present after fabrication or developing during service. The degree of

attention which the engineering industry has focussed on each of these three

types of material has varied from time to time, depending on the major current

problems.

In the early days of brittle fracture, the major cause of concern was the

fracture toughness of the parent material, leading to the wide spread use of

the Charpy impact test to classify materials, as an aid to selection of steel

for applications of varying severity. Other studies of the factors affecting

toughness, led to a long process of improvements in steel composition and

processing, has led to a situation where, it is possible to conclude that, for

the majority of structural applications, the range of steels currently

available is adequate to guarantee safety against fast fracture over those

parts of the structure unaffected by welding.

As steel properties have improved, and as welded structures have developed into

larger and more complex designs, increasingly operating under more severe

conditions and with serious commercial and safety penalties in the event of

failure,

complementary developments in welding consumables and procedures have

been necessary. The offshore oil and gas field exploitation, and the

accompanying distribution systems and processing plants, have resulted in a

marked increase in fundamental work and commercial development work on both

manual and automatic consumables and procedures, such that the properties

required, particularly fracture toughness, have been achievable in massive

complex structures. Defect tolerance approaches to design, facilitated by

improved stress calculation methods and inspection techniques, have played a

vital role.

Interest in the HAZ has varied from time to time, in technical terms, and has

in more recent years assumed strong commercial aspects.

1.2.1 Technical Importance of HAZ Properties

In the past, the importance of HAZ toughness has been closely considered with

the incidence of hydrogen induced cold cracking, on which a vast amount of

research work has, and is being carried out. It is now clearly

understoodt

1

-^ )

that three conditions must be present for H2 cracking to

occur :-

An HAZ with a susceptible hardened microstructure,

promoted by higher carbon equivalent values (CEV)

and faster cooling rates.

A weld metal hydrogen content above a critical

level for the given steel and type of joint.


34/204

FR 62-10 822 7210.KA/804

A critical level of stress across the joint,

varying depending on the severity of the two

factors,

given above, promoted by heavy

restraint in the joint, and intensified by

any unfavourable geometric features present.

Based on existing data, monograms are available

(

l

"

5

) which for the normal

range of structural grades, will allow hydrogen cracking to be avoided by weld

metal hydrogen control and/or adjustment of cooling rate by preheating to a

temperature which is defined by the CEV and combined thickness.

Investigational work continues into such areas as joints of greater thickness,

the effect of high yield stress (in, for example, Q & Τ steels), and the

behaviour of the latest types of structural steel (e.g. low carbon, very low

sulphur).

Until relatively recently, the properties of the HAZ have only been regarded in

the sense of their ability or otherwise to tolerate the presence of spasmodic

H2 cracking. The relative defect tolerance of the HAZ against other potential

failure modes has, by comparison, received much less attention, historically,

perhaps for the very good reason that it is almost impossible to discover any

failures in which the HAZ has been the prime cause, except in the presence of

H2 cracks.

The measures taken to design out H2 cracking (lowering of CEV, reduced cooling

rates, etc.) will also tend to improve the fracture toughness of the HAZ - a

dual gain in joint reliability. Therefore, the potential failure modes are

initiation from lack of fusion/penetration defects, after fabrication, or due

to environmental effects during service (fatigue, stress corrosion, corrosion

fatigue).

The toughness required to withstand pre-existing fabrication defects can be

assessed using well established defect tolerance approaches, based in this case

on the grain coarsened HAZ properties, together with a knowledge of the service

stresses and NDT discrimination level.

The situation with respect to in-service cracking is extremely complex,

particularly when corrosive environments are involved. Design procedures

against fatigue in normal atmospheres are well establishedÍ

1

-°) and now being

included in British Standards. HAZ's are of interest because the likely

initiation sites at the edges of the weld reinforcement, for instance, will

ensure that the growing crack will, firstly, have its tip located in the HAZ

for the first part of its growth, and, secondly, will, during this period, be

subjected to the maximum stress concentration factors (SCF) due to the weld

profile and any geometric effects associated with the weld location. It is

relatively easy, however, to define a worst case after the potential fatigue

initiation sites have been identified and to calculate the fracture toughness

required to ensure the survival of the structure over this interval of fatigue

crack growth.

A number of factors combine to reduce the probability of failure from fatigue

cracking:-

The region of lowest toughness, the grain coarsened

HAZ,

usually has the highest strength, and an advancing

crack will tend to turn rapidly away into softer

material, of higher toughness. This phenomenon

presents constant problems in the fatigue precracking

of HAZ test pieces, particularly where cracks adjacent

to the fusion boundary are required. The outer areas

of the HAZ and the parent plate, which come into play

as the crack grows, are of higher toughness and can provide

the necessary increased defect tolerance, once this

initial danger period is over.

In many instances, e.g. fillet welds, the SCF associated

with the fatigue crack initiation sites decays very

rapidly, and this effect coupled with the probability

of increasing crack tip toughness, can cause a progressive

decrease in crack growth rate and even lead to stable


35/204

FR 62-10 822 7210.KA/804

arrested cracks, which will be below the maximum

tolerable size for the crack tip material and, hence,

present no danger to the integrity of the structure.

In corrosive atmospheres, particularly for steels in contact with media which

charge the steel with hydrogen, the situation may be less encouraging.

Hydrogen has the general effect of reducing the resistance of steels to almost

any form of cracking, so that two factors, at least, can lead to an increased

possibility of failure:-

The crack tip fracture toughness, per se, may be

reduced, resulting in a lower tolerable defect size.

Crack growth can continue under lower stress ranges,

making the situation of sessile non-critical cracks

less likely.

The relatively high strength of some parts of the HAZ, in comparison to the

parent plate, and the presence of residual stresses from welding can sensitise

the areas adjacent to welds to certain forms of environment cracking. In some

cases,

the resistance of the parent steel to more generalised types of

cracking will be important.

The consideration of these environmental effects, which are of prime interest

at present in the oil and gas recovery and processing fields, are, however,

outside the scope of this report, but may have some influence on attitudes

towards the properties of welded joints under less severe conditions. Steel

and/or weld procedure selection may be constrained, for instance, by a maximum

HAZ hardness criterion.

Another important aspect which has reduced the effect of local embrittlement,

is the narrowness of the band of material worst affected, generally 1 mm or

less. Unless the properties are at or, near the lower shelf of the transition

behaviour, or loading is extremely rapid, the plastic zone at the crack tip

will be able to grow, under increased loading, to such an extent that tough

material on either side becomes involved in the prefracture process zone,

giving increased crack tip toughness. Ductile crack growth, if it occurs and

the mechanical constraints on crack propagation are not too severe, will tend

to favour movement into softer and tougher HAZ regions. There is a close

similarity to the situation referred to above for advancing fatigue cracks, in

that once an initial danger period is past, the situation at the crack tip can

improve markedly. This inability of most HAZ's to contain a critical event is

the most likely reason for the absence of failure case histories citing the

properties of sound HAZ's as a cause of failure, and has led to the

recommendation in the UK that it is unnecessary to carry out HAZ testing, on

the normal types of structural steels, where the weld energy lies between 1.5

kJ/mm and 4-4.5

kJ/mmt

1

-

7

)-

There is, of course, great commercial interest in

increasing productivity by using higher weld energies, but it becomes advisable

then to assess the effect of HAZ degradation, since the increased HAZ width may

allow the HAZ to assume a potentially more significant role in possible failure

modes.

Heat affected zones in electroslag welds are an extreme case, being not

only extremely wide, of the order of 5-8 mm, but having an extended region of

very coarse austenitic grain size adjacent to the fusion boundary. The

combination of low toughness and large width make the reliability of the HAZ

for structural purposes very suspect. The weld metal, similarly, with a coarse

as-cast structure, has very poor toughness, and it has been usual to normalise

electroslag welds intended for service at ambient temperatures. The economic

advantages of the process still motivate development work, directed towards

improving the as-welded properites of electroslag welds, such as narrow gap

welding to reduce heat input and improve economicsÍ

1

-

8

), improved alloyed weld

metals

I

1

-

9

),

and the use of steels more tolerant of high heat inputs(

x

-

1

°).


36/204

FR

62-10 822

7210.KA/804

Temperature, C

1500

1400.

1300-

1200-

1100

ifr

ìsai

7 Effect of varying

interpass temperature

Time, s

HEAT AFFECTED ZONE THERMAL CYCLES - 5 kJ/mm HEAT INPUT

FIG. 1.1


37/204

F R 6 2 - 1 0 8 2 2

7 2 1 0 . K A / 8 0 4

WELD MACROGRAPH (5 kJ/mm, SINGLE V)

SHOWING REHEATED HEAT AFFECTED ZONES

FIG.

1.2


38/204

FR 62-10 822 7210.

KA/804

2. EXPERIMENTAL PROGRAMME

2.1 Aims and Obj ect ives of the Work

In view of the foregoing discussion of the increasing requirements of customers

and classification authorities for HAZ testing, and the difficulties which the

inhomogeneous and narrow nature of the HAZ presents in testing, there were two

discrete but complementary major aspects to the programme:-

(a) A critical examin ation of the available methods for

investigating the fracture properties of the HAZ's

of structural stee ls, with particular reference to

their suitability for guaranteeing fitness for purpose

of welded s truc ture s, with the overall aim of assess ing

the relationships, if any, between sub-size and full

thickness te sts, and ascertaining the degree of safety

of defect tolerance predictions, using wide plate tests.

(b) Using current high quality structural steels of various

types, to gauge the practical effect of the above tests

in,

so far as possible, real situations within the

present fabrication industry.

To achieve these aims, three steels in current production were chosen:-

Fe510DD (European

25-72),

aluminium grain refined

but containing no niobium.

BS4360:Gra de 50D, an aluminium killed, niobium

grain refined steel of equivalent minimum

guaranteed tensile and impact toughness properties.

RQT 500, a BSC proprietary quenched and tempered

steel,

of similar composition to BS4360:50D,

but having higher gu arant eed t ensile and impact

energy lev els (470 N/m m

2

YS, 41 J at -40°C).

Alth ough the majori ty of the work was carried out on 40 mm thick pl at es ,

additional plates at 25 and 60 mm were included to define thickness effects

arising from thermal differences in the welds and varying constraints in full

thickness test pieces.

To preserve the relevance to current practice, tests were extracted from full

thick ness welds with various ty pes of weld prepar atio n (single and double V, K,

butt),

using prac tica l weld pro ce du re s, over a wide range of heat input (2 and

5 kJ/mm plus electroslag), to gauge weld energy and HAZ width effects.

As well as the effec ts re sulting from the metho dolog y of tes ting , it was of

interest to attempt to gain information on the underlying reasons for

differences in behaviour between weld types and heat inputs and between

different steel compo sitio ns. All welds were examined metallog raphical ly and

small scale tests were also produced by a thermal simulation technique using

specially built equip ment , programmed with thermal data derived from the test

welds.

Details of the various aspects of experimental work will be given in subsequent

sections.

2.2 Steel s Used in the Test Prog ramme

Types of steels commonly used in high demand applications, such as offshore

structures,

were chosen, over a range of composition types and yield strengths.

No attempt was made to use other than normal commercial steels; full width

plates were ordered through normal chan nels . Full ultrasonic examination was

requested to avoid wast age of test materi al due to internal de fe ct s. The

Fe510DD plate ex Creusot Loire was ultrasonically examined to NFA 04 305,

normal grade, and the BS4360:50D and RQT 500 plates, both of BSC manufacture,

to BS5996 Grade LC3.


39/204

FR 62-10 822 7210.KA/80 4

The chemical compositions and mechanical properties of all the steel plates

tested are given in Tables 2.1 and 2.2 , respectiv ely. Results are given from

the mill certificates and check tests carried out on the plates received.

2.2.1 Normali sed Steels

Both BS43 60:5 0D and Euron orm 25/72 : Fe510DD are steels aimed at the same

market,

high demand appli cations for welded fabrications in high yield ste el,

requiring low temperature impact energy guar ante es. Minimum property levels

are very similar in the two st ee ls . The major diff eren ce between the two

lies in the approach to grain refinement; BS4360:50D uses niobium additions to

an aluminium killed CMn steel, Fe510DD relies on the presence of A1N to

achieve grain refinement - there is no niobium additio n. The two philosophi es

have implications related to welding, with respect to the effects of the

relative CEV's required to achieve the guaranteed tensile properties and to the

effic iency of the grain refining mechanis m in the coarse grained HA Z. Both of

these have potential for affecting HAZ properties.

Compa ring the two 40 mm sample plat es in Table s 2.1 and 2.2 , it will be

that:-

(a) The chemical analyses are satisfa ctory, except that

the % C in two of the Fe510DD sa mples was margi nally

over the maximum 0.22% allowed in the product, which

together with a slightly higher Mn content and the

presence of some Cr, gave a CEV of 0.51, compared

with 0.40 in the 50D st ee l. The margin over the

minimum guaranteed yield strengths was higher in

Fe5 10D D. On the other hand , the 50D sample

achieved better impact toughness levels , at -45°C ,

than the Fe510DD at -20°C, in both testing

directions.

BS4360 :50D was used to compare thickness effe cts, and plates at 25 , 40 and 60

mm thick were teste d. Tabl es 2.1 and 2.2 show that the comp osit ions wer e

similar, the major differences lying in the carbon contents, and hence

CEV's.

All plates met the minimum tensile requi remen ts, with rather small margins in

the 40 and 60 mm plates, but a rather large margin of -40 N/mm

2

in the 25 mm

plates,

no doubt because of its rather high C conte nt. Neve rthe less , its

CEV was only 0.42%. All samples wer e, simila rly, comfortably above the

minimum guaranteed impact toughness levels, the 25 and 60 mm plates being

similar, at a much lower level than the 40 mm plate, bearing in mind the test

temperature difference.

The microstructures observed in the plates received are shown in Fig. 2.1.

They are essentially all of the same type - bands of pearlite in a ferritic

matrix, but there are differences, apparent even at the optical microscopy

level. The higher carbon conten t of the 40 mm Fe510DD plate is observed as a

higher volume fraction of pear lite . Centreline segregation is obvious in the

Fe510DD and 50D samples, causing bainitic bands to appear.

2.2.2 Quench ed and Tempe red Stee l

In order to look at the effects of chemical composition, initial microstructure

and strength level on HAZ struct ures and prop ert ies , it was decided to include

as one of the sample mater ial s, RQT 500, a BSC proprietary roller-quenched and

tempered steel . These steels are produced from, basicall y, the same feedstock

as BS4360:50D, the higher property levels being achieved by virtue of the

difference in heat treatme nt.

In Table 2.1, it will be seen that, comparing the 40 mm plates R4 and D4, the

CEV's are very similar, the lower carbon level in D4 being offset by its higher

Mn, Cu and Ni con ten ts. The 25 mm pla te, R2 , is lower in CEV, almost totally

due to lower carbon content.

Table 2.2 shows, firstly, that the plates achieve the minimum guarantee levels

comfor tably with one marg inal yield level in R4 . The longi tudina l impact

energies are comparable with those obtained in the normalised D4 plate,


40/204

FR 62-10 822 7210. KA/80 4

although greater anisotropy is obvious from the transverse resu lts. Second ly,

the QT treatment has raised the yield strength and UTS by -130 N/mm

2

.

The micro stru ctur es of the two RQT 500 plate s are shown in Fig. 2.2, and can be

seen to consist of equiaxed ferrite, with bainitic areas.

2.3 Weld Procedur es

The main body of the work was based on large test we ld s, aimed at :-

(a) Allowing the practic al difficulti es of the

various types of test to be assessed.

(b) Indicat ing the prope rty levels achieved by

the three steel grades used, over a range

of test types, weld c onfigurations and

energy in puts, and plate thic kness es.

So far as pos sib le, the weld proced ures laid down used normal fabric atio n sh op

methods,

to avoid creating a laboratory type study which would not be of

general applicability in the fabrication industry.

Welds were carried out on the common types of weld preparation - single V,

double V and K, plus square butt preparati ons for the electroslag we lds . For

the 25 mm plat es on ly, the K format was modified to a single bev el (90° one

side,

45° on the other), Fig. 2.6. The programme aimed to cover, also , as

wide a range of weld energy inputs as possible, from manual metal arc (MMA)

levels, through submerged arc (SA) heat in puts , up to typical elec trosl ag (ES)

levels.

Beca use of the inherent variabil ity in heat inputs in MMA we ld s, it

was decided to use two submerged arc machine settings, at 2 kJ/mm, typical of

MMA , and 5 kJ/mm, a typical input energy for single wire SA welding and

slightly above the upper limits of the range where HAZ properties of normal

structural steels are widely regarded as being of no practical concern í

1

-

7

).

The energy input to ES welds is around an order of magnitude higher, depending

on operating cond itio ns. In the 40 mm pla tes , for example, it was around 50

kJ/mm , but is diff icult to compare because of conducti ve losses to the copper

cooling shoes.

Preheat levels were applied according to the recommendations of B S S l S o t

1

-

5

) , on

the basis of combined joint th ick nes s, CEV and heat inp ut, to avoid HAZ

hydrogen induced cracking, which could, of course, have interfered seriously

with the testing pr ogr amm e. In order to avoid pote ntia l prob lems due to

fatigue cr acks , correctly located in the HA Z, breaking across into the weld

metal, as has occurred in some HAZ studies(

2

-

) ,

extremely tough SA weld metals

were prod uced using a Mo-B wire (Oerlikon Tibor 22) with a fully basic fl ux,

O P 1 2 1 T T (

2

-

2

).

Heavy strong backs were used to restrain the two halve s of the

test weld, to simulate a real structure and prevent, so far as possible,

angular distor tion, which can cause problems in COD testing. In cutting back

the original root runs prior to welding the second side of a weld , a minim um

amount of grinding was used, particularly in the Κ welds, to prevent

dist urba nce to the flat HA Z. No air-arc gouging was used becaus e of the HAZ

which it can pro du ce. This is not unduly seve re, but it was diff icul t to

assess the potential heat input, and because of its location at the weld root

might affect the heat input effects which the test programme was intended to

evaluate.

Except for the weld preparations and heat inputs, the general methods used were

identical and not all weld proc edur es will be det aile d. Exampl es are however

given in Figs. 2.3-2.6.

The large number of test specimens required necessitated the production of

2.5 m long we ld s. For con sis ten cy, these were carried out as a single weld ,

with the exception of the electroslag wel ds, where the machine size limitations

necessi tated the prep arat ion of two 1.5 m we ld s. After wel din g, the plat es

were examined ultrason ically for major defe cts, which might have caused

specimen wastage.

10


41/204

FR 62-10 822 7210.KA/804

2.4 Test Methods

Whenever possible, standard test methods and/or practices used in the

fabrication industry were employed, although in many instances no specific

instructions are given for the HAZ situation.

2.4.1 Charpy V-Notch Impact Tests

Charpy impact energy tests were carried out on 10 χ 10 mm standard specimens,

to BS131:Part 2. Check tests on sample plates were carried out in longitudinal

and transverse orientations, in the normal subsurface positions. In addition,

for comparison with root specimens in the two-sided welds, specimens in the

transverse orientation were extracted from the 40 and 60 mm plates. Specimens

were tested over a temperature interval to define the transition behaviour.

The specimens taken from the weld HAZ's, in subsurface and root locations

(Fig. 3.9) were marked at the fusion line (FL) and at 1 mm, 2 mm and in some

cases 5 mm into the HAZ. All specimens were transverse to the rolling

direction, to take the worst case, and the notches were oriented along the

rolling direction (TL specimen). Ten specimens were tested for each notch

location, to define the transition curve, for comparison with the parent plate

values.

In order to locate the notches, overlength specimens were etched in the finish

machined (10 χ 10 mm) size and the notch position marked by scribing. The

notches were then machined and the specimen trimmed to length. In all cases,

the FL reference was taken at the centre of the face of the specimen. In the

welds with sloping fusion boundaries, single and double V welds, the notch in

the surface specimen will therefore include about 50% weld metal in FL

specimens and progressively less at the other notch locations. Depending on

the exact shape of the fusion boundary, the root specimens and the specimens

taken from Κ and ES welds should have less weld metal involved in the notch

root.

2.4.2 Sub-size (10 χ 10 mm) Crack Opening Displacement Tests

Specimens for this test were extracted in identical fashion to the Charpy V

test specimens - subsurface and root locations, TL orientation. Specimens

were again macroetched to locate the notch positions, notching in this case

being carried out with a 0.15 mm thick rubber bonded SiC cutting wheel,

followed by fatigue extension. The provisions of BS5762:1979(*-

3

) were

followed with respect to notch geometry, testing and the calculation of crack

tip COD from the test results. Testing was carried out on a screwdriven

Instron machine (150 kN

capacity),

with a cold N2 gas atmosphere, maintained by

a feedback control system, at the test temperature, by means of a thermocouple

attached to the specimen.

2.4.3 Full Thickness COD Tests

These were carried out to the full provisions of BS5762:1979 on 2B χ Β TL

oriented specimens. The specimens were again macroetched in machined blank

condition and notch positions marked as before, for cutting with a V tipped

diamond cutter, prior to fatigue extension. Four specimens were normally used

for each notch location, with the aim of defining, if possible, the transition

range.

After testing, the broken test pieces were sectioned at right angles to the

fracture face, across the position of fracture origin, if visible, or at the

centre of the thickness, if not, polished and examined microscopically to

determine the position of the fatigue crack tip relative to the fusion

boundary.

Tests were carried out on an Avery-Denison hydraulic machine, of 50 t capacity.

2.4.4 Wide Plate Tests

Wide plate tests were carried out on a 4000 t facility, the test plate, 1 m

square,

was welded, after the longitudinal test weld has been notched, into two

11


42/204

FR 62-10 822 7210.KA/804

loading yokes against which the four 1000 t rams bear. Cooling was carried out

with liquid-nitrogen sprays, over the central area of the plate.

The notches in the six 5 kJ/mm K welds tested (40 mm thick, two from each

grade,

at -30 and 40°C) were surface notches with a nominal 10:1 aspect

ratio, so far as the deepest part of the notch was concerned. A 6 mm deep

pre-notch was cut into the weld so that its tip lay in the HAZ, nominally 1 mm

from the fusion boundary. Judging the position on the surface of the weld,

ground flush with the plate surface, was extremely difficult, since the only

points of reference were the ground and macroetched ends of the weld, some

0.5 m from the area of interest, and the fusion boundary of the capping bead on

the surface in the area to be notched. Any changes in the cross section of

the weld, or the width of the capping bead along the length of the weld, were

likely to lead to lateral misplacement of the notch tip. Mechanical

prenotching was carried out with a specially designed rotary saw, with a

0.15 mm SiC blade, which could be moved along the length of the notch to

produce a flat tip, with ends following the radius of the blade. This prenotch

was subsequently extended, by bending fatigue, to about 9 mm in total. The

actual values are given in Table 3.10. The progress of fatigue extension was

followed using a precalibrated ac impedance device.

The plate was instrumented using linear displacement transducers (LDT) and COD

clip gauges as shown in Fig. 2.7. A range of gauge lengths was used for the

LDT's to show, to an extent, the strain distribution in the plates. The

strain on each LDT was sampled at approximately 5 s intervals during the test

and stored, with the instantaneous load reading, in a data logger. The three

clip gauges were mounted in the centre and at the two extremities of the notch,

and their outputs were logged in the same fashion. In addition, normal COD

test XY plots against load were produced. After the test, the fracture

surfaces were photographed, and then sectioned at right angles, as in the COD

tests,

to determine the notch position relative to the fusion boundary.

2.4.5 Thermal Cycle Data Acquisition

The interest in the thermal history stemmed from a desire (a) to know the

general levels of heating rate and cooling rate over the approximate

transformation range, taking in common with many other studies as 800-500°C

(Ats-s), in order to produce representative HAZ material by thermal simulation,

and (b) to observe any systematic differences in cooling rate between the

different weld preparation types used. Details of the method used to embed

thermocouples close to the fusion face of a representative selection of the

actual test weld plates are presented in the Appendix, together with the

results obtained.

2.4.6 Simulation Studies

A simulator was built, for this programme, using resistance heating, from a

welding transformer, of 12 χ 12 mm transverse specimens, held in water cooled

jaws, which were capable of imposing cooling rates in excess of those recorded

from the test welds. Thermal cycles typical of various regions in the HAZ's

of the test welds were imposed using a Research Inc. DATA-TRAK unit controlled

by feedback from a Pt-Pt/Rh thermocouple, spot welded to the centre of the heat

treated region of the specimen.

The histograms given in the Appendix suggest that the most characteristic Ate-5

values for 2 kJ/mm, 5 kJ/mm and electroslag welds, in 40 mm plate, are 15, 48

and 245-295 s respectively. Simulated HAZ specimens were produced for a single

pass situation, i.e. no tempering effects, for peak temperatures of 950, 1150

and 1350°C, with Ate-5 values of 16 and 50 s, to cover 2 and 5 kJ/mm welds,

and, for electroslag welds, a single peak temperature, 1300°C and a Ate-5 of

300 s.

To gain some insight into the effects of tempering by subsequent runs,

specimens from each steel were given an initial thermal cycle to 1350°C and asecond to 950°C, the Ate-s value in each case being 50 s.

Charpy V impact toughness tests were carried out in each simulated condition,

together with Vickers hardness tests at 5 kg load (HV5) and optical microscopic

examination.

12


43/204


44/204

TABLE 2.2 MECHANICAL PROPERTIES OF PLATES TESTED

TI

Steel Grade

Fe510DD

(Normalised)

BS4360:50D

(Normalised)

Thickness

mm

40

(DD4)

25

(D2)

40

(D4)

60

(D6)

Mill Cert.

Check 1

Check 2

Check 3

Spec.

Mill Cert.

Check 1

Check 2

Check 3

Spec.

Mill Cert.

Check 1

Check 2

Check 3

Spec.

Mill Cert.

Check 1

Check 2

Check 3

Spec.

YS

N/mm

2

(L)405

(L)384

(T)388

(L)398

(T)400

(L)404

(T)405

>345

(L)392

(L)397

(T)398

(L)397

(T)391

(L)391

(T)396

>355

(L)426

(L)365

(T)351

(L)347

(T)355

(L)353

(T)362

>345

370

(L)367

(T)357

(L)361

(T)358

(L)342

(T)342

>340

UTS

N/mm

2

585

574

572

583

588

583

590

510/610

540

542

550

546

550

549

547

490/620

525

511

499

503

505

506

504

490/620

545

526

525

516

523

519

512

490/620

Elong,

on

5.65 / S

% °

29

40

32

37

37

39

>22

29

28*

31

28

26

30

31

>20

32

43

42

41

44

43

41

>20

32

31*

32

33

39

30

32

20

Impact Energy, J

(L)

(T)

(T)

(T)

(L)

(L)

(T)

(T)

(T)

(L)

(T)

(T)

(T)

(T)

(L)

(L)

(T)

(T)

(T)

(L)

96,

72, 102 at -20°C

76,

84, 84; (L) 124, 114, 120

77,78,

82; (L) 94, 103, 11

afc

_

2()0(;

64, 81, 84; (L) 109, 126, 81

40 J average at -20°C

67, 57, 64 at -30°C

78,

100; (L) 72, 90

64, 80; (L) 103, 90

afc

_

2 Q O C

84,

76; (L) 78, 87

41 J average at -20°C, 27 J average at -30°C

57, 68, 59; (L) 94, 103, 128 at -40°C

140, 140, 150; (L) 90, 126, 136

102,

139, 128; (L) 116, 74, 121

afc

_

i50Q

134, 96, 126; (L) 130, 139, 128


88,

116, 98 at -20°C

98, 88; (L) 96, 108

78, 70; (L) 120, 96

afc

_

2QO(:

96,

89; (L) 92, 84



45/204

TABLE 2.2

(Continued)

τι

55

Steel Grade

RQT 500

(Quenched

and

Tempered)

Thickness

mm

25

(R2)

40

(R4)

Mill Cert.

Check 1

Check 2

Check 3

Spec.

Mill Cert.

Check 1

Check 2

Check 3

Spec.

YS

N /mm

2

545

(L)359

(T)528

(L)506

T)

524

(LJ519

(T)520

>470

564

(L)501

(T)502

(L)485

(T)469

(L) -

(T)490

>470

UTS

N/mm

2

645

636

636

620

630

624

625

560/710

680

625

629

619

611

638

560/710

Elong,

on

5.65

fS

% °

46

42

40

42

40

42

40

>21

24

20

19

19

24

22

>21

Impact Energy, J

(L) 115, 130, 108 at -40°C

(T) 86, 86, 74; (L) 130, 126, 124

(T) 80, 76, 78; (L) 110, 105, 115 at -40°C

(L) 41 J average at -40°C

(L) 62, 66, 80 at -40°

(T) 91, 81, 85; (L) 92, 118, 82

(T) 75, 87, 94; (L) 106, 104, 86

afc

_

4 Q O C

(T) 53, 86, 7 8; (L) 178, 120, 84

L 41 J average at -40°C

Converted from non-standard gauge length to a gauge length

of 5.65 /S according to BS3894:Part 1, 1965

>


46/204

FR 62-10 822 7210.KA/804

TABLE 2.3

SUB-DIVISION OF WELD TEST PLATE

^ A Β C

E

G

A

Test Piece Allocation

mm

A -

Β -

C -

D -

E ■

F ■

G ■

Discard

- 2B χ Β Full thickness COD

Reserve

Macro/micro/analysis

Thermometry

Charpy impact toughness

10 χ 10 mm COD

80

1145

255

50

205

560

205

2500

(R1/8872)

16


47/204

FR 62-10 822

7210.KA/804

^«^^a^:


48/204

FR 62-10 822

7210.KA/804

x 200

Subsurface

χ 400

Subsurface

Centre

(b) BS4 360 .50 D, 25 mm thick

FIG. 2.1

Continued)


49/204

F R 6 2 - 1 0 822 7 2 1 0 . K A / 8 0 4

Tr-· ^ ^ ^ * * * * * * ^

: ^ ^ > r ^ '

ψ0* *'Μ

^ ' Ρ * ' - V .

x 200 Subsurface

x

400 Subsurface

ι*»*̂ ^ *̂***̂

1

?**̂ *̂ ̂ .,. i > (j*-

m** y

v

^ ■ ^ ^ ^ ^ ^ ^ Τ feüfät V i w *&&

x 200 Centre

x

400 Centre

(c) BS4360:50D, 40

mm

thick FIG.

2.1

(Continued)

19


50/204


51/204

F R 62- 1 0 822

7 2 1 0 . K A / 8 0 4

χ

200

S u b s u r f a c e

mmmmmm

. Ä ^ V ' - . * - » « * · .

1

' Ο

· . -

Τ, V ' / * ' · ; · · ·' '.

V- . ' . ' · \

Vãt?r-v-?-»s*v

l

-^-V'j ' . . . ' ' . . 'ν ¿--■'«'ν·''·.

<

f.

'■■-

ν . ; Γ Λ · . ·- '·' "

iv . -ΐν'· ·: ·., ■··' -Γ **" '.·-■ ,-.ν . ..·

'S.¿\:

•·*^-··':

ν

r

' i r

-Vi.·.

. ■ · ' · > ■ ·ν . ·· .· ■· > ■ - « · > . , ' · . ·· · - y .

X

200

C e n t r e

χ 400

C e n t r e

M I C R O S T R U C T U R E S OF A S - R E C E I V E D Q U EN C HE D AND T E M P E R E D P L A T E S

RQT

5 00 , 25 mm

T H I C K

F I G . 2. 2

2 1


52/204

F R 6 2 - 1 0 8 2 2

7210 .KA/804

WELD No.

D4K5

(W673)

PROJECT

ECSC

HAZ

PROGRAMME

NUMBER

7210 .KA/804

SIGNED A .

Welder

DATE

WELD PREPARATION

S e e over

PASS tOCATION AND SEQUENCE

S e e over

PARENT

PLATE

THICK

(mm)

4 0

GRADE

STOCK No

B S 4 3 6 0 : 5 0 D

WELDING PARAMETERS

W 3

POWER SUPPLY

S I D E 1

d c

SIDE 2

d c

METHOD OF PREPN.

Flame c u t

POLARITY

+ v e

+ v e

WELDING PROCESS

Submerged a r c

POSITION

Downhand

ELECTRODE

CONFIGURATION

S i n g l e

S i n g l e

TOET

1

ONSUMABLES

TYPE

STORE

No.

PASS No.

4 - 1 2

FILLER WIRE

R o o t

-

E7016-1

R p s t

-

T i h n r

? ?

DIAMETER mm)

HOT WIRE

STICK OUT mm)

3 0

Documents

Measurement of toughness in the heat affected zone of welded structural steels